fracture characterization of pc/abs blends: effect of ... · 23 700 8500 7.7 0.39 mitsubishi...

Fracture characterization of PC/ABS blends: effect of reactivecompatibilization, ABS type and rubber concentration

G. Wildes, H. Keskkula, D.R. Paul*

Department of Chemical Engineering and Center for Polymer Research, The University of Texas at Austin, Austin, TX 78712, USA

Received 5 January 1998; accepted 12 August 1998

Abstract

The fracture of thin (3.18 mm) and thick (6.35 mm) specimens of polycarbonate (PC), acrylonitrile–butadiene–styrene (ABS) and PC/ABS blends with both standard and sharp notches was examined by standard Izod and single edge notch three point bend (SEN3PB)instrumented Dynatup tests. Ligament lengths were varied for SEN3PB samples and the corresponding fracture data were represented byplotting the specific fracture energy (U/A) as a function of ligament length (`). The slope of these plots was found to be closely related to boththe Izod impact strength and the size of the process (or stress whitened) zones surrounding the crack surface. Significant morphologycoarsening was seen in some PC/ABS (70/30) blends for thick (6.35 mm) injection molded parts; blends compatibilized with 1% of an SAN-amine polymer exhibited well-dispersed, stable morphologies. The effects of blend composition, ABS type, sample geometry, PC molecularweight, temperature, rubber concentration and reactive compatibilization on the fracture properties of these materials were examined.q 1999 Elsevier Science Ltd. All rights reserved.

Keywords: Polycarbonate; Acrylonitrile–butadiene–styrene; Acrylate–styrene–acrylonitrile

1. Introduction

The commercial success of blends of polycarbonate (PC)with acrylonitrile–butadiene–styrene (ABS) materials iscaused by a combination of factors such as improvedprocessability relative to PC and high fracture toughnessin the presence of sharp cracks, in thick sections and atlow temperatures [1–15]. It is useful to characterize thefracture performance of PC/ABS blends in some detail.The theory of linear elastic fracture mechanics (LEFM)restricts analysis to materials with only small-scale yieldingnear the crack-tip. Because the plastic zone size is signifi-cantly larger than the size of test specimens in ductile engi-neering polymers, measurement of meaningfulK1c values isoften not realistic for such materials. Traditionally, the frac-ture performance of engineering polymer blends has beencharacterized by the Charpy or Izod tests. Although Izodtesting with a single ligament size offers a simple andconvenient method for comparing many materials, the frac-ture characterization of toughened polymers with varyingligament sizes offers a more thorough description of thecrack formation and propagation behavior of these

materials. Many recent articles have shown the benefits ofanalysis of fracture data using a two parameter model, whichanalyzes the fracture energy per unit area as a function of theligament size [10,16–26]. The use of this type of analysishas provided effective characterization of toughened poly-mer blends in previous reports from this laboratory [27,28].Use of thick samples with sharp notches under impact load-ing conditions offers severe test conditions which betterdifferentiate between super tough polymeric materials thatmay have very similar Izod impact values. This articleexamines the effects of reactive compatibilization, blendcomposition, sample geometry, temperature, and rubberconcentration on the fracture properties of PC/ABS blends.

2. Analysis of fracture data with varying ligament size

Two mathematical approaches for analyzing fracturebehavior as a function of ligament size were introduced.Each is based on the idea first introduced by Broberg thatthe region at the tip of a crack consists of an end regionwhere actual fracture occurs and an outer region whereenergy is plastically absorbed during crack propagation[29].

The essential work of fracture (EWF) method is based onthe partitioning of the fracture energy of ductile materials

Polymer 40 (1999) 7089–7107

0032-3861/99/$ - see front matterq 1999 Elsevier Science Ltd. All rights reserved.PII: S0032-3861(98)00865-9

* Corresponding author. Tel.:1 1-512-471-5392; fax: 1 1-512-471-0542.

E-mail address:[email protected] (D.R. Paul)

into two components, the essential and non-EWF, asproposed by Mai and coworkers [16,21,30,31]. The essen-tial work (We) is related to the formation and localizeddeformation associated with crack extension, while thenon-essential work term (Wp) incorporates the plasticwork in the volume surrounding the crack region. Thesum of these two terms gives the total work of fracture (Wf).

Wf We 1 Wp: 1For a given sample thickness,t, We is proportional to the

ligament length,̀ andWp has a volumetric dependence andis therefore proportional tò 2

Wf `twe 1 `2tbwp; 2whereb is the plastic zone volumetric shape factor. Thespecific EWF (energy per unit area) iswe We/`t whilewp Wp/`2tb is the specific non-EWF (energy per unitvolume) equivalent to the energy dissipated in the plasticzone. Therefore, a normalized expression for the totalspecific work of fracture,wf, is as follows:

wf we 1 bwp`: 3Thus, a plot ofwf versus̀ should give a linear relationship

where the intercept at zero ligament length iswe and theslope isbwp. For PC,we has been experimentally shown tobe independent of notch geometry, specimen width, andspecimen length [17].

The fracture energy,U, of ductile materials with varyingligament sizes has been proposed by Vu Khanh [24–26] tofollow a relationship of the form

U GiA 1 1=2TaA2; 4whereA `t is the area of the ligament. The intercept of aplot of U/A versusA or Gi was called the fracture energy atcrack initiation while the termTa related to the slope wascalled a tearing modulus or the rate of change ofGr (definedasGi 1 TaA) with crack extension.

In addition to the differences in interpretation of the slopeand intercepts of these similar linear plots, there are differ-ences in the method of loading, sample geometry, and test-ing rate in the approaches used by Mai et al. and Vu Khanh.Table 1 compares typical testing parameters and samplegeometries from the literature for the two methodologies;sample geometry is schematically shown in Fig. 1. It shouldbe emphasized that while there are similarities in the math-ematical analysis of the data obtained from these tests, theactual test conditions are very different; the EWF methoduses very thin specimens which are tested at low rates inuniaxial tension, while Vu-Khanh’s methodology employshigh speed loading of relatively thick specimens in bendingwhich induces a triaxial stress state in the material. Clearlythe fracture characterization of toughened polymers withvarying ligament sizes offers a more thorough descriptionof the crack formation and propagation behavior of thesematerials compared to Izod testing which uses a singleligament size.

Many recent articles have shown the benefits of analysisof fracture data using a two-parameter model

G. Wildes et al. / Polymer 40 (1999) 7089–71077090

Table 1Test conditions and sample geometry for various fracture methods

Test Test method Test velocity Stress condition t (mm) W (mm) Z (mm) Ligament̀ W 2 a (mm)

EWF SENT, DENT-Tensile (razornotch)

1.7 × 1025 m/s Plane stress 1.6 25 (25–50) 58 (20–70) 5–45; (0.2–0.9)W

Izod Cantilever-impact (standardnotch)

3.5 m/s Plane stress–strain 3.2 12.7 62 10.2; 0.8W

Vu Khanh SEN3PB-impact (razor notch) 2.5, 3.0 m/s Plane stress–strain 3–6 10, 12 50, 90 2–10; (0.2–0.8)WThis study SEN3PB–impact (razor notch) 3.5 m/s Plane stress–strain 3.2, 6.4 12.7 5.4 2–10; (0.2–0.8)W

Fig. 1. Schematic of test specimen geometry; values used in this study andin the essential work of fracture, Izod, and Vu Khanh methods are given inTable 1.

Fig. 2. Schematic of the reaction of SAN-amine with bisphenol-A-polycarbonate to form SAN-g-PC[42].

G. Wildes et al. / Polymer 40 (1999) 7089–7107 7091

Tab

le2

PC

and

SA

Nm

ater

ials

used

inth

isst

udy

Pol

ymer

Des

igna

tion

Des

crip

tion

� Mw

� Mn

Tor

que

(Nm

)aR

elat

ive

mel

tvi

scos

ityS

ourc

e

Cal

ibe

r20

0-3

VH

-PC

BP

Apo

lyca

rbon

ate

Ver

yhi

ghM

W36

000

1340

016

.80.

17D

owC

hem

ical

Com

pany

Iupi

lon-

E20

00H

-PC

BP

Apo

lyca

rbon

ate

Hig

hM

W32

000

1080

013

.10.

23M

itsub

ishi

Eng

inee

ring

Pla

stic

sC

orpo

ratio

nIu

pilo

n-S

3000

M-P

CB

PA

poly

carb

onat

eM

ediu

mM

W23

700

8500

7.7

0.39

Mits

ubis

hiE

ngin

eeri

ngP

last

ics

Cor

pora

tion

Iupi

lon-

H30

00L-

PC

BP

Apo

lyca

rbon

ate

Low

MW

2010

074

004.

60.

65M

itsub

ishi

Eng

inee

ring

Pla

stic

sC

orpo

ratio

nLu

stra

n35

SA

N32

.5S

tyre

neac

rylo

nitr

ileco

poly

mer

,32

.5%

AN

130

000

5900

03.

001.

00B

ayer

Cor

pora

tion

Tyr

il10

00S

AN

25S

tyre

neac

rylo

nitr

ileco

poly

mer

,25

%A

N15

200

077

000

3.6

0.83

Dow

Che

mic

alC

ompa

ny

SA

N-a

min

eS

AN

-am

ine

Am

ine

func

tiona

lS/A

N/a

min

e(6

7/32

/1)

117

300

4810

03.

80.

79B

ayer

Cor

pora

tion

aB

rabe

nder

torq

ueta

ken

afte

r10

min

mix

ing

at27

08C

and

60rp

m.

[10,17,18,21,26–28,30]. This study uses thick specimenswith sharp notches at high rates of loading in order to betterdiscern between super tough PC/ABS blends and elucidatethe importance of constituent materials and morphologieson the fracture performance of these materials. As will beshown, the specific fracture energy of the materials used inthis study is more closely a function of ligament length thanarea. By plotting the normalized fracture energy as a func-tion of ligament length, it is recognized that the interceptand slope may not be equal towe andbwp since the testconditions and sample geometries employed do not conformexactly to the yielding criterion required in the EWF analy-sis. Therefore, we have defined the intercept,u0 ;lim`!0U=A; simply as the limiting specific fractureenergy, and the slope,ud d(U/A)/d`, as the dissipative

energy density in the following relationship

U=A u0 1 ud`; 5where ud has units of energy/volume and it presumablyreflects plastic deformation energy in the process (or stresswhitened) zone surrounding the fracture surface. Thisanalysis method is introduced as an adaptation of theEWF concept to facilitate the performance evaluation ofsuper tough blends with thick parts and at high speeds.

3. Compatibilization of PC/ABS using SAN-amine

Reactive compatibilization of polymer blends by the insitu formation of block or graft copolymers reduces


Table 3ABS and ASA materials used in this study

Polymer Designation Description Soluble SAN Torque (N m)a Source

�Mw �Mn

Magnum541 ABS16 16% Rubber 140 000 59 000 3.9 Dow Chemical Company25% AN in SAN

Lustran38 ABS38 38% Rubber 130 000 59 000 10.0 Bayer Corporation30% AN in SAN

SAN-g ABS45 45% Rubber 90 000 35 000 12.7 Cheil Industries25% AN in SAN

BL65 ABS50 50% Rubber 167 000 44 000 12.8 Sumitomo NaugatuckCorporation

24% AN in SANGeloy ASA46 46% Acrylate rubber NA 9.6 GE Plastics

33% AN in SAN

a Brabender torque taken after 10 min mixing at 2708C and 60 rpm.

Fig. 3. Brabender torque versus time for H-PC, ABS38, and H-PC/ABS38 (70/30) blends with and without SAN-amine compatibilizer mixed at 2708C and arotor speed of 60 rpm (16% rubber in ABS phase).

interfacial tension, introduces a steric hindrance to coales-cence and can improve the interfacial adhesion betweenphases. Chemical schemes for the compatibilization ofpolyamide blends using maleated rubber modifiers arewell known [32–41]; however, the absence of functionalchain ends on PC does not provide a straightforward routefor compatibilization of its blends. The formation of SAN-g-PC copolymer at the PC/SAN interface is accomplishedthrough the chemical scheme shown in Fig. 2 [42]. TheSAN-amine polymer is synthesized by reacting a styrene/acrylonitrile/maleic anhydride (67/32/1) terpolymer with1-(2-aminoethyl) piperazine in a reactive processingscheme. The SAN-amine polymer is miscible with theSAN matrix of ABS and reacts with PC as shown. Thegraft copolymer formed was shown to reduce the SANdispersed phase particle size for different blend composi-tions, AN contents of the SAN, PC molecular weights andprocessing conditions [42–44]. The purpose here is toexamine the effects of reactive compatibilization on themorphology dependent fracture properties of PC/ABSblends.

4. Experimental

4.1. Materials

Details of the PC and SAN polymers employed in thisarticle are shown in Table 2. The designation for eachcommercial PC is based on its relative molecular weight;e.g., very high molecular weight polycarbonate (VH-PC).Commercial SAN copolymers containing 32.5% AN(SAN32.5) and 25% AN (SAN25) were used; the function-alized SAN-amine polymer was described in a previousarticle [42]. Table 3 describes the characteristics and suppli-ers of the ABS and the acrylate–styrene–acrylonitrile(ASA) materials used in this article; each is designated bythe type of material followed by its as-received rubber phaseconcentration. The ABS and ASA materials employed inthis article represent different rubber concentrations(16%–45%), AN contents (24%–32%), rubber types(acrylate- and butadiene-based elastomers), and manufac-turing processes (bulk and emulsion polymerization). Allmaterials were dried in a vacuum oven at 808C for at least12 h before melt processing. Fig. 3 shows Brabender torqueas a function of time for H-PC, ABS38 diluted to 16%rubber, and 70/30 blends of these materials with and without1% SAN-amine. As seen in Fig. 3, the melt viscosity of PC/ABS blends can be significantly lower than for PC which isan important processing advantage of these materials.

Blends were prepared by melt mixing at 40 rpm in a Kill-ion single screw extruder (L/D 30, 25.4 mm diameter)


Fig. 4. Schematic of Dynatup single edge notch three point bend (SEN3PB) experimental configuration (10 kg tup at a test velocity of 3.5 m/s) [28].

Fig. 5. Schematic of injection molded fracture bar indicating gate and farend fracture specimens; thicknesses of 3.18 mm and 6.35 mm were used.

outfitted with an intensive mixing head. Low rubber contentABS materials were made by blending virgin ABS withSAN25 or SAN32.5, corresponding to the AN content ofthe ABS material to be diluted. The ABS pellets of reducedrubber concentration were then molded into test specimensor blended with PC. PC/ABS blends were prepared byblending PC with either a virgin ABS or a diluted ABS.ABS materials were processed at 2008C while PC/ABSblends were processed at 2708C. There was no significantdifference in the morphology or properties of blendsprepared by pre-blending compared with those mixed in aone step process. Test specimens for mechanical propertyevaluation were produced using a 75 ton Arburg Allrounder305 screw injection molding machine.

Izod impact testing was performed using a TMI 15 Jpendulum type tester. The impact strength data representaverage values of at least six test specimens. Standarddeviations for these data are typicallŷ 5%, but the

variability is greater near transition regions. Standardnotch test bars (thickness 3.18 mm) conformed toASTM D 256. Liquid carbon dioxide was used to coolsamples for low temperature impact testing and a thermo-couple implanted in a neighboring test specimen was used tomonitor the temperature.

4.2. SEN3PB methodology

Instrumented impact tests were made using a DynatupDrop Tower Model 8200 with a 10 kg weight at a fracturevelocity of 3.5 m/s as shown in Fig. 4 [28]. The single notch,three-point bend (SEN3PB) specimen geometry is describedin Table 1 and Fig. 1. Between 24 and 36 samples (one halfgate-end and one half far-end specimens as shown in Fig. 5)were tested with original ligament lengths between 2 and10 mm, (0.2–0.8)W. A sharp notch was made by pressing afresh razor blade cooled in liquid nitrogen into the root of


Fig. 6. Fracture energy as a function of ligament area for (a) ABS45 (45% rubber) and (b) H-PC/ABS38 (70/30) with 16% rubber in the ABS phase for thin(3.18 mm) and thick (6.35 mm) samples with a sharp notch.

the notch. Fracture energy was calculated from the inte-grated area under the load-deflection curves. AlthoughIzod impact testing of these super tough materials resultedin partial breaks with significant unbroken ligaments,Dynatup testing of SEN3PB samples led to complete breaksor hinged ligaments that were generally 0.25 mm or smallerin length. The test apparatus and methodology are describedin more detail in previous articles from this research group[27,28].

4.3. TEM analysis

Transmission electron microscopy (TEM) was used toexamine the morphology of the melt processed blends ofPC/ABS and PC/ABS/SAN-amine. Injection molded speci-mens were microtomed with a diamond knife perpendicularto the plane of flow in the center of the test sample. Thecryogenically microtomed sections, about 20 nm thick,

were prepared using a Reichert-Jung Ultracut E at a sampletemperature of2108C using a knife temperature of2458C.

The butadiene rubber phase of the ABS was stained(black in TEM photomicrographs) by exposing microtomedsections to the vapor of a 2.0% aqueous solution of OsO4 atroom temperature for 18 h. The PC phase of the blends waspreferentially stained (gray in TEM photomicrographs) byexposing the ultra-thin sections to the vapors of a 0.5%aqueous solution of RuO4 at room temperature for 9 minafter staining with OsO4. The dual staining techniqueprovides effective contrast between all three phases in theblend: PC, SAN and rubber. Staining with RuO4 beforeOsO4 resulted in poor phase contrast; the presence ofRuO4 seems to reduce the effectiveness of OsO4 as a stain-ing agent. Blends of PC with ASA were stained in a one stepprocess which provided contrast between all three phases byexposing ultra-thin sections to the vapors of a 0.5% aqueoussolution of RuO4 at room temperature for 8 min. The use of


Fig. 7. Fracture energy as a function of ligament length for (a) ABS45 (45% rubber) and (b) H-PC/ABS38 (70/30) with 16% rubber in the ABS phase for thin(3.18 mm) and thick (6.35 mm) samples with a sharp notch.

RuO4 to stain acrylate-based rubber has been reported else-where in the literature [45,46]. TEM imaging was done on aJeol 200CX microscope operating at 120 keV.

5. Fracture evaluation of PC/ABS blends

5.1. Effect of sample thickness

Fig. 6(a) and (b) show the fracture response as a functionof ligament area,A `t, for ABS45 and H-PC/ABS38 (70/30) blends at two different sample thickness. These twoplots suggest that the interceptu0 is relatively independentof sample thickness for the test methodology and specimengeometry used in this article. However, as seen in Fig. 6(a),the slope is decreased by a factor of three when the speci-men thickness is doubled. Figs. 7(a) and (b) show that plotsof fracture energy versus ligament length,`W2 a, ratherthan area result in a slope that is relatively independent ofsample thickness compared to plots of fracture energy as afunction of ligament area. This suggests that the fractureenergy for these materials and testing conditions is moretruly a function of ligament length than ligament area.

5.2. Effect of PC molecular weight

The fracture characteristics of the PC materials used inthis article are given in Table 4. Increased PC molecularweight results in higher impact strength for 3.18 mm and6.35 mm specimens with both standard and sharp notches.Thin (3.18 mm) samples with a standard notch are supertough at room temperature while all samples with a sharpnotch have Izod impact strengths less than 100 J/m. Thickspecimens (6.35 mm) failed in a brittle manner for bothstandard and sharp notch samples; samples with standardnotches exhibited Izod impact strengths approximatelyfive times higher than those with sharp notches. The fracturesurface of thick specimens showed that samples with stan-dard notches had a larger rough area near the original cracktip; the greater surface roughness could account for thehigher impact energies observed. The limiting specific frac-ture energy,u0, and the dissipative energy density,ud, werecalculated by plottingU/A as a function of ligament lengthfor specimens 6.35 mm thick. All four PC materials resultedin plots of U/A versus` with zero slope, i.e.ud 0; u0ranged from 4.5 for the highest molecular weight to2.5 kJ/m2 for the lowest molecular weight PC.

The values ofu0 obtained here are similar to relatedparameters reported in other studies; Paton and Hashemi[17] reported a plane-strain specific work of fracture of3 kJ/m2 and a specific work of fracture initiation of 4.3 kJ/m2 using the EWF method. Plati and Williams [47], andlater Singh and Parihar [48], reported values of 4.8 and6.13 kJ/m2, respectively using theJ-integral technique.Each of these testing methods characterizes the plane-strainfracture response of materials; therefore, calculations weredone to confirm that the tests described in this article were


Tab

le4

Fra

ctur

epr

oper

ties

ofpo

lyca

rbon

ate

mat

eria

ls

3.18

mm

Spe

cim

ens

6.35

mm

Spe

cim

ens

Sta

ndar

dno

tch

Sha

rpno

tch

Sta

ndar

dno

tch

Sha

rpno

tch

PC

Izod

impa

ctst

reng

th(J

/m)

Duc

tile

–br

ittle

Izod

impa

ctst

reng

th(J

/m)

Izod

impa

ctS

tren

gth

(J/m

)Iz

odim

pact

stre

ngth

(J/m

)D

uctil

e–

britt

leP

lane

Str

ain

frac

ture

para

met

ers

tran

sitio

ntr

ansi

tion

tem

pera

ture

(8C)

tem

pera

ture

(8C)

u 0(K

J/m

2 )u d

(MJ/

m3 )

VH

-PC

1000

242

9524

550

.60

4.5

0H

-PC

980

240

9022

545

.60

4.3

0M

-PC

880

230

8520

040

.60

2.5

0L-

PC

790

210

7514

530

.60

2.5

0

also done under plane-strain conditions. The load versustime data from the Dynatup was used to estimate the strainrate during testing, approximately 150 s21; the effectiveyield stresss y 79 MPa at this rate was calculated usingan Eyring plot for PC [49]. Using a published value of thestress intensity factor ofK1c 2.24 MPa m1/2 from Ref.[50], the minimum specimen thickness defined by

tmin 2:5K1c=sy2 6for plane-strain fracture conditions as recommended byASTM is calculated to be approximately 2 mm. Thus, theuse of the 6.35 mm specimen thickness assures plane-strainfracture for PC as used in this article.

Fig. 8 shows the Izod impact strength of PC/ABS16(70/30) blends with three different PCs as a function of

temperature using thin (3.18 mm) samples with a standardnotch. As seen in Fig. 8, higher PC molecular weights leadto higher impact strengths and lower ductile–brittle transi-tion temperatures in these blends. Although the impactstrength of the neat PC materials was higher than the corre-sponding PC/ABS16 (70/30) blend for thin (3.18 mm)samples, all PCs were brittle in thick (6.35 mm) sectionswhile the blends weretough. H-PC was chosen for allfurther blend studies because of its combination of lowtemperature toughness and lower melt viscosity comparedto VH-PC.

5.3. Effect of ABS material on PC/ABS blend properties

Table 5 shows the impact fracture characteristics of the


Fig. 8. Izod impact strength of PC/ABS16 (70/30) blends as a function of temperature for VH-PC, H-PC and M-PC using thin (3.18 mm) samples with astandard notch.

Table 5Fracture property comparison of ABS and ASA materials

3.18 mm 6.35 mm

Standard notch Standard notch Sharp notch

Plane strain fracture parameters

Polymer Izod impact strength (J/m) Izod impact strength (J/m) u0 (KJ/m2) ud (MJ/m

3)

ABS16 340 235 22 0ABS38 540 270 15.5 1.0ABS45 550 340 10.0 2.7ASA46 540 340 16.6 1.9ABS50 520 340 10.2 2.6ABS16a 235 22 0ABS38a 65 6.8 0ABS45a 80 8.6 0ASA46a 85 8.9 0

a Properties of the ABS or ASA materials indicated after blending with appropriate SAN copolymers to obtain a final rubber content of 16% by weight.


Fig. 9. Izod impact strength of H-PC/ABS (70/30) blends for thin (3.18 mm) samples with a standard notch.

Table 6Fracture properties of H-PC/ABS38 blends with 16% rubber in the ABS phase

3.18 mm 6.35 mm


H-Pc/ABS38 blend Ductile–brittle Izod impact Strength (J/m) Ductile–brittle Izod impact Strength (J/m) Plane strain fracture parameterstransition transitiontemperature (8C) temperature (8C) u0 (KJ/m

2) ud (MJ/m3)

100/0 240 980 .60 225 4.3 070/30 224 785 10 675 18.7 4.450/50 218 775 210 440 19.5 3.40/100 250 55 65 6.8 0

Fig. 10. Dynatup fracture energy for H-PC/ABS38 blends as a function of ligament length for thick (6.35 mm) samples with a sharp notch.


Fig. 11. (a) Izod impact strength, (b) limiting specific fracture energy (u0), and (c) dissipative energy density (ud) as a function of blend composition for H-PC/ABS blends with 16% rubber in the ABS phase for thick (6.35 mm) samples.


Fig. 12. Dynatup fracture energy (U/A) of H-PC, ABS38 and H-PC/ABS38 blends with 16% rubber in the ABS phase for thick (6.35 mm) samples with a sharpnotch and a ligament length of 10 mm.

Table 7Comparison of PC/ABS and PC/ASA (70/30) blends with 16% rubber in the ABS phase

3.18 mm 6.35 mm


Blend (70/30) Ductile–brittle Izod impact strength (J/m) Izod impact strength (J/m) Plane strain fracture parameters(16% rubber transitionin ABS) temperature (8C) u0 (kJ/m

2) ud (MJ/m3)

H-PC/ABS16 223 770 535 11.2 5.5H-PC/ABS38 224 785 675 18.7 4.4H-PC/ABS38(11%SAN-amine)

226 775 670 18.9 4.4

H-PC/ABS45 229 790 725 16.7 5.5H-PC/ASA46 224 900 745 13.6 5.1

Fig. 13. Ductile–brittle transition temperature as a function of rubbercontent in the ABS phase for H-PC/ABS38 (70/30) blends using thin(3.18 mm) samples with a standard notch.

Fig. 14. Izod impact strength as a function of rubber content in the ABSphase for H-PC/ABS38 (70/30) blends using thin (3.18 mm) and thick(6.35 mm) samples a standard notch.

different ABS and ASA materials used in this study. Aspreviously discussed, each material has a different rubberconcentration and morphology. For comparison, two sets offracture properties are reported for these materials, the as-received materials and these materials diluted with the appro-priate SAN to a fixed (16%) rubber concentration. As might beexpected, the materials with high rubber contents have some-what higher Izod impact strength values for both 3.18 and6.35 mm specimens. The mass produced ABS16 has the high-estu0 while the high rubber content emulsion materials havehigher values ofud. At a fixed rubber content of 16%, onlyABS16 shows high values of impact strength for thick samples(6.35 mm) and was the only material to exhibit any stresswhitening near the fracture zone.


Fig. 15. Fracture energy as a function of ligament length for H-PC/ABS38 (70/30) blends with varying rubber concentrations in the ABS phase for thick(6.35 mm) samples with a sharp notch.

Fig. 16. Dissipative energy density (ud) of far-end and gate-end specimens (6.35 mm thick, sharp notch) as a function of rubber content in the ABS phase for H-PC/ABS38 (70/30) blends using thick (6.35 mm) samples with a sharp notch.

Fig. 17. Schematic of SEN3PB fracture specimen indicating outer,S andinner,S0, stress whitened zone sizes measured after cryo-fracturing one halfof the sample.

Izod impact strength of blends of H-PC with ABS16,ABS38, and ABS 45 using thin (3.18 mm) samples areshown in Fig. 9. While all three blends have similar roomtemperature impact strength, the blends with higher rubbercontent ABS materials have superior low temperaturetoughness compared to the blend with ABS16; the blendof H-PC/ABS45 (70/30) was ductile to2708C which wasthe limit of the CO2 cooling apparatus used in this study.ABS50 was not used for blending with PC because it hasbeen previously shown to severely degrade PC at processingtemperature [51].

5.4. Effect of blend composition

The fracture characteristics of H-PC, ABS38 and H-PC/ABS38 blends are shown in Fig. 10 for thick samples

(6.35 mm) with sharp notches. Blends of H-PC with 30%and 50% ABS38 exhibited superior toughness at all liga-ment lengths compared to unblended H-PC and ABS38.However, to effectively compare the extent of tougheningprovided by different ABS and ASA materials, each mate-rial was diluted with SAN to give the same rubber concen-tration before blending with H-PC. The effect of blendcomposition on the fracture properties of blends of H-PCwith ABS16, ABS38, ABS45 and ASA46 at a fixed rubbercontent of 16% in the ABS phase are compared in Fig.11(a)–(c). Fig. 11(a) shows the thick (6.35 mm) specimenIzod impact strength as a function of ABS content in H-PCblends. While 70/30 blends of H-PC with ABS16, ABS45,ABS38, and ASA46 are all super tough, blends with ABS45,ABS38, and ASA46 have superior fracture properties forblends with 50% PC. As seen in Fig. 11(c), very toughblends with corresponding high values ofud can beproduced from H-PC and ABS materials that are brittle orsemi-brittle under the test conditions used in this study andhaveud 0. Izod values seem to be more dependent on theslope, ud, than the intercept,u0, of U/A versus` plots.However,u0 does clearly affect impact strength as seen bycomparing the various ABS materials shown at 0% PC inFig. 11(a)–(c).

Addition of ABS38 to H-PC reduced the room tempera-ture Izod impact strength and increased the ductile–brittletransition temperature for thin (3.18 mm) samples, butsignificantly increased the Izod impact strength, decreasedthe ductile–brittle transition temperature and increasedbothu0 andud for thick (6.35 mm) samples. Fig. 12 clearlyshows the synergistic effect of blending H-PC with ABS38containing 16% rubber in the ABS phase. H-PC/ABS38blends exhibited greater toughness at all temperaturescompared to the unblended materials using thick(6.35 mm) samples. A more detailed analysis of the frac-ture properties of H-PC/ABS38 blends with a fixed rubber


Fig. 18. Outer stress whitened zone size,S, as a function of rubber contentin the ABS phase for H-PC/ABS38 (70/30) blends using thick (6.35 mm)samples with standard and sharp notches.

Fig. 19. Dissipative energy density (ud) versus. outer stress whitened zone size,S, using thick (6.35 mm) samples with a sharp notch.

content of 16% in the ABS phase is summarized in Table6.

5.5. Effect of ABS type on the fracture properties of PC/ABS(70/30) blends

Table 7 compares the fracture behavior of H-PC blendswith different ASA and ABS materials at a fixed blendcomposition of 70% PC and with a total rubber concentra-tion of 4.8%. To produce a mixture with this rubber concen-tration, blends of H-PC/ASA and H-PC/ABS with 70%H-PC were diluted with SAN to give a rubber concentrationof 16% rubber in the ABS and ASA phase. At this fixedblend composition and rubber content, the emulsion-madeABS and ASA materials exhibit superior toughness whenblended with PC. Blends of H-PC with ABS45 at this lowtotal rubber concentration (4.8%) had both the lowestductile–brittle transition temperature and the highest valueof ud. Blends of ABS16 with H-PC are also very tough;however, this mass-made ABS leads to lower fractureenergies for thick specimens compared to blends ofABS38, ABS45, and ASA46 with H-PC. The use ofbutyl acrylate rubber instead of butadiene rubber inASA materials significantly improves weatherability butmight be expected to result in inferior impact propertiesunder certain conditions because of the higherTg of therubber phase. However, the blend of H-PC with ASA46exhibited the highest impact strength for thin (3.18 mm)samples, similar ductile–brittle transition temperatures asblends of H-PC with ABS, and excellent thick specimenfracture characteristics.

5.6. Effect of rubber content in ABS

The rubber concentration in the ABS phase of H-PC/ABS38 (70/30) blends was varied by pre-blending ABS38

with SAN32.5. Fig. 13 shows that for thin samples with astandard notch increasing the rubber concentration in theABS phase above the critical rubber concentration (10%)gradually decreases the ductile–brittle transition tempera-ture of the blend. Decreasing the rubber concentrationbelow 10% in the rubber phase significantly increases theductile–brittle transition temperature to over 608C for H-PC/SAN32.5 (0% rubber).

Fig. 14 shows that for room temperature toughness thereis a critical rubber concentration in the ABS phase. For thick(6.35 mm) samples with a standard notch, this criticalrubber concentration was found to be 10% rubber in theABS phase (3% total rubber in the blend), while thin(3.18 mm) samples with a standard notch were tough evenat 5% rubber in the ABS. At the critical rubber concentra-tion for thick samples, the gate end of the specimen exhib-ited tough behavior with significant stress whitening but thefar end of the specimen was only semi-tough and showedareas of brittle fracture. The ductile–brittle behavior forthick specimens is also shown in Fig. 15 by plottingDynatup fracture energy as a function of ligament length.The slope,ud, of theU/A versus̀ plot for H-PC/ABS38 (70/30) as a function of rubber concentration is shown in Fig.16; this parameter seems closely related to Izod impactstrength asud is zero until the critical rubber content isreached where there is a step change in value followed bya gradual increase with rubber concentration.

Many tough polymeric materials exhibit a stress whitenedzone, which surrounds the fracture surface. The size of thisstress whitened zone can be measured perpendicular to thefracture surface; however, the extent of stress whitening inthe interior of thick specimens is not necessarily the same asseen on the surface of the sample because of volumetricconstraints on the material during loading. In thick speci-mens like those used here, plane strain conditions tend to


Fig. 20. Dissipative energy density (ud) as a function of rubber content of the ABS phase for H-PC/ABS38 (70/30) blends with 1% SAN-amine compatibilizerusing thick (6.35 mm) samples with a sharp notch.

exist in the interior of the sample while the surface experi-ences plane stress conditions. The relationship between theouter (S) and inner (S0) size of the stress whitened zone isschematically shown in Fig. 17 where one half of theSEN3PB specimen was fractured after cooling in liquidnitrogen to expose the cross-sectional area of the sample.Fig. 18 shows the outer stress whitened zone size (S) as afunction of rubber concentration for both standard and sharpnotch thick (6.35 mm) samples. The size of the stresswhitened zone was measured as the maximum width ofthe zone perpendicular to the crack on the outer surface ofthe specimen as shown in Fig. 17. Fig. 18 shows a similardependence on rubber content as that seen in Figs. 15 and 16with the same critical rubber concentration (10%). Above10% rubber, the stress whitened zone size increases almostlinearly with rubber concentration, while the impactstrength andud are relatively independent of rubber content

in this range. These seemingly contrary observations can beexplained by the fact that while the outer stress whitenedzone size (S) increases, the inner size of the zone (S0) isnearly constant for rubber concentrations greater than10%. Therefore, both the fracture energy and the totalvolume of stress whitened material change only nominallyabove the critical rubber concentration. The relationshipbetween these parameters is shown in Fig. 19 with a lineardependence ofud on the outer stress whitened zone size ofthick (6.35 mm) Izod specimens.

5.7. Effect of reactive compatibilization

Comparison of Figs. 20 and 16 shows that incorporationof 1% SAN-amine in a 70/30 blend of H-PC/ABS38 signif-icantly improves the fracture toughness of the far end of theinjection molded sample. The difference in fracture


Fig. 21. Fracture energy as a function of ligament length for (a) H-PC/ABS38 (70/30) and (b) H-PC/ABS38 (70/30) using 1% SAN-amine compatibilizer, with10% rubber in the ABS phase for thick (6.35 mm) samples with a sharp notch.

properties between the far and gate ends of thick (6.35 mm)samples for blends with 10% rubber in the ABS phase isseen in Fig. 21(a) and (b). This difference was thought to bethe result of morphological rearrangements in the moldduring the filling and cooling cycles. The effect of reactivecompatibilization using the previously described SAN-amine polymer on the morphology development andstability of PC/SAN was examined in two previous articles[43,44]. Although compatibilized and uncompatibilizedblends exhibited similar morphologies and properties forthin (3.18 mm) samples, there were significant differencesfor thick injection molded specimens. Figs. 22(a)–(d) showthe dramatic difference between the morphology of the farends of the injection molded 6.35 mm samples for compa-tibilized and uncompatibilized H-PC/ABS38 at the criticalrubber concentration (10% in the ABS phase). The ABS

phase was well dispersed in the thin samples and the gate-end of the thick samples as shown in Fig. 22(b). Thissuggests that ABS can be well dispersed in PC withoutthe aid of a compatibilizer, but the poor dispersion seen inFig. 22(a) results from the morphological instability ofuncompatibilized PC/ABS blends. The morphology of thecompatibilized blend is much more stable.

6. Conclusions

The fracture of thin (3.18 mm) and thick (6.35 mm)specimens of PC, ABS and PC/ABS blends with both stan-dard and sharp notches was examined by standard Izod andsingle edge notch three point bend (SEN3PB) instrumentedDynatup tests. Ligament lengths were varied for SEN3PB


Fig. 22. TEM photomicrographs showing the morphology of the (a) far-end and (b) gate-end of injection molded bars for uncompatibilized H-PC/ABS38 (70/30) blends with 10% rubber in the ABS compared to (c) far end and (d) gate-end of H-PC/ABS38 (70/30) blends with 10% rubber in the ABS compatibilizedusing 1% SAN-amine.

samples and the corresponding fracture data were repre-sented by plotting the specific fracture energy (U/A) as afunction of ligament length (̀). Blends of PC with ABS oracrylate–styrene–acrylonitrile (ASA) were found to besuper tough for thick and thin samples with both standardand razor notches. These blends had higher fracture energyvalues at zero ligament length, intercept u0, and dissipa-tive energy densities, slope ud, compared to either PC orABS materials. Neat PC materials with higher molecularweight have larger intercepts but they remain independentof ligament length, i.e., the slope orud remains 0. For PC/ABS (70/30) blends, a critical rubber concentration of 10%in the ABS phase (3% total rubber in the blend) was foundwhere the far-end of thick (6.35 mm) injection molded barswere brittle (ud 0) and the gate-end was tough (ud . 0);ud increased with increasing rubber content above this criti-cal value.

The incorporation of 1% of an SAN-amine compatibili-zer produced tough samples for both far and gate-endsamples at the same rubber concentration. TEM analysisof the blend morphology confirmed that the loweredfracture toughness in the far-end of the uncompatibilizedspecimens was because of the formation of large ABSdomains caused by coalescence during the injection mold-ing cycle; both the gate and far-end of the compatibilizedblend had small, well dispersed ABS domains. Izod impactenergy andud were shown to be directly proportional to thestress whitened zone size surrounding the fracture surfaceof the samples. The presence of a stress whitened zoneseems to be a necessary but not sufficient requirement fora material to exhibit a significantud; there were someblends with a stress whitened zone butud 0, whilethere were no blends with a positiveud without a stresswhitened zone.


Fig. 22. (continued)

Acknowledgements

This research was supported by the US Army ResearchOffice. Various polymers used in this work were supplied byBayer Corporation, Cheil Industries, The Dow ChemicalCompany, GE Plastics, Mitsubishi Engineering PlasticsCorporation and Sumitomo Chemical. The authors aregrateful to Clive Bucknall for his helpful discussions andinsight. The authors would also like to thank Allen Padwa ofBayer Corporation for his contributions in this work.

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